5. Overvoltages And Insulation Co-ordination: Internal Origin

390

5. OVERVOLTAGES AND INSULATION CO-ORDINATION Different types of overvoltage may occur in industrial networks. Devices must therefore be installed to reduce their magnitude and the insulation level of equipment must be chosen so that fault risks are reduced to an acceptable level.

5.1.

Overvoltages

An overvoltage is any voltage between one phase conductor and earth, or between phase conductors having a peak value exceeding the corresponding peak of the highest voltage for equipment, defined in standard IEC 71-1. An overvoltage is said to be of differential mode if it occurs between phase conductors or between different circuits. It is said to be of common mode if it occurs between one phase conductor and the frame or earth.

n origin of overvoltages Overvoltages can be of internal or external origin.

o internal origin

These overvoltages are caused by a given network element and only depend on the characteristics and structure of the network itself. For example, the overvoltage that occurs when a transformer's magnetizing current is interrupted.

o external origin These overvoltages are caused or transmitted by elements outside the network, for example: - overvoltage caused by lightning - spread of HV overvoltage through a transformer to the internal network of a factory.

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n classification of overvoltages Standard IEC 71-1 gives the classification of overvoltages according to their duration and form. According to the duration, a distinction is made between temporary overvoltages and transient overvoltages: - temporary overvoltage: power frequency overvoltages of relatively long duration (from several periods to several seconds). - transient overvoltage: short-duration overvoltage lasting only several milliseconds, which may be oscillatory and is generally highly damped. Transient overvoltages are divided into: . slow-front overvoltage . fast-front overvoltage . very-fast-front overvoltage.

n standard voltage forms Standard IEC 71-1 gives the standardised wave forms used to carry out tests on equipment: - short-duration power frequency voltage: this is a sinusoidal voltage with a frequency between 48 Hz and 62 Hz and a duration equal to 60 s. - switching impulse: this is an impulse voltage having a time to peak of 250 µs and a time to half-value of 2500 µs. - lightning impulse: this is an impulse voltage having a front time of 1.2 µs and a time to half-value of 50 µs. n consequences of overvoltages Overvoltages in electrical networks cause equipment degradation, a drop in service continuity and are a hazard to the safety of persons. The consequences can be very varied depending on the type of overvoltages, their magnitude and their duration. They are summed up as follows: - breakdown in the insulating dielectric of equipment in the case where the overvoltage exceeds the specified withstand - degradation of equipment through ageing, caused by non-destructive but repetitive overvoltages - loss of power supply caused by the destruction of network elements

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- disturbance of control, monitoring and communication circuits by conduction or electromagnetic radiation - electrodynamic stress (destruction or deformation of equipment) and thermal stress (elements melting, fire, explosion) essentially caused by lightning impulses - hazard to man and animals following rises in potential and occurrence of step and touch voltages. 5.1.2.

Power frequency overvoltages

Power frequency overvoltages are generally caused by: - an earth fault - resonance or ferro-resonance - neutral conductor breakdown - a generator voltage regulator or transformer on-load tap changer fault - overcompensation of reactive energy following a varmeter regulator fault - load shedding, notably when the supply source is a generator 5.1.2.1.

Overvoltage caused by an earth fault

Overvoltages caused by the occurrence of an earth fault greatly depend on the neutral earthing system of the given network.

n unearthed (MV or LV) or impedance earthed (MV) neutral Figure 5-1 shows that on occurrence of a solid earth fault, the voltage between the neutral point and earth becomes equal to the single-phase voltage: VNeutral = Vn Vn : nominal single-phase voltage For a fault on phase 1, VNeutral = − V1 . The phase-earth voltage of healthy phases thus becomes equal to the phase-to-phase voltage: V2 E = VNeutral + V2 = V2 − V1 V3E = VNeutral + V3 = V3 − V1 whence

V2 E = V3E = 3 Vn

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V3 3

V2

2

V1 1

VNeutral

ZNeutral

V2

earth fault

Neutral

V1E

V2E

V3E

V3

V3E

V2E V1

V1E

0

V1 , V2 , V3 : phase-neutral voltages V1E , V2 E , V3E : phase-earth voltages Z Neutral : earthing impedance ( Z Neutral = ∞ for an unearthed neutral) Figure 5-1: overvoltage on an unearthed or impedance earthed network on occurrence of a phase-to-earth fault Note 1 : for an impedance earthed neutral, the value of Z Neutral is much greater than the value of the transformer and cable impedances and the fault resistance, which is why VNeutral = − V1 . Note 2 : in overhead public distribution networks, there are highly resistive faults (several kΩ), having a value close to or higher than the earthing impedance. In this case, a highly resistive fault will cause an overvoltage lower than 3 Vn .

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n solidly earthed neutral (HV or MV) On occurrence of an earth fault on one network phase, a high current is generated which circulates in the circuit formed by the fault phase, earth and neutral earth electrode (see fig. 5-2). At the fault point, the three-phase voltage system is disturbed. The fault phase voltage in relation to earth is almost zero if we neglect the fault resistance. The voltages of the other two phases in relation to earth are higher than the single-phase voltage, while remaining lower than the phase-to-phase voltage. V3 ZT

ZC

ZT

ZC

ZT

ZC

V2 V1

fault

V1E

V3E

Rf

Re

V1 , V2 , V3 ZT ZC Re Rf

V2 E

: single-phase voltages : transformer impedance : cable impedance : neutral earth electrode resistance : fault resistance

Figure 5-2: equivalent diagram of a phase-earth fault when the neutral is solidly earthed

Thus, we can define an earth fault factor k characterising the phase-earth overvoltage occurring on the healthy phases: V2 E = V3E = k Vn Vn : nominal single-phase voltage The symmetrical component calculation method (see § 4.2.2. of the Protection guide) can be used to determine the value of k in relation to the positive, negative and zero-sequence impedances:

k = 1−

Z(1) + a 2 Z( 2 ) + a Z( 0 ) Z(1) + Z( 2) + Z( 0 ) + 3 R f

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In most networks, generators are sufficiently far away to take the approximation Z(1) = Z( 2 ) ; we thus have:

k = 1+

(

a Z(1) − Z( 0)

)

2 Z(1) + Z(0 ) + 3 R f

Nomographs can be used to determine factor k for a zero fault resistance ( R f = 0 ) in relation R(0) X(0 ) and for R(1) = 0 and R(1) = 0.5 X(1) (see fig. 5.3. et 5.4.). to the ratios X(1) X(1) where: R(1) : positive-sequence resistance seen from the fault point X(1) : positive-sequence reactance seen from the fault point R(0 ) : zero-sequence resistance seen from the fault point X( 0) : zero-sequence reactance seen from the fault point When the fault resistance is not zero, we can see in the formula expressing k that the overvoltage is weaker. The calculation of the overvoltage with a zero fault resistance thus provides an excess value. If we again use the diagram in figure 5-2, we can determine these impedances for a practical case: by taking:

ZT = RT + j XT ZC = RC + j XC

  positive - sequence impedances 

Z( 0 )T = RT + j X( 0 )T    zero - sequence impedances Z( 0 )C = RC + j X( 0 )C 

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we can determine: R(1) = RT + RC X(1) = XT + XC R(0) = 3 Re + RT + RC X(0 ) = X(0 )T + X(0)C Note:

A factor 3 appears before Re . The reason for this is explained in figure 4-11 of the Industrial network protection guide. R(0) 8 X (1) 7

k = 1.7 6

k = 1.6

5 4 3

k = 1.5 2

k = 1.4

1

k = 1.3

k = 1.2 1

2

3

4

5

Figure 5-3: earth fault factor in relation to ratios

6

X(0 )

7

and

X(1)

8

X(0) X (1)

8

X(0) X (1)

R(0) X(1)

for R(1) = 0 and R f = 0 R(0) 8 X (1) 7

k = 1.7 k = 1.6

6

k = 1.5

5 4

k = 1.4 3 2

k = 1.5

k = 1.3 k = 1.2

1

1

2

3

4

5

Figure 5-4: earth fault factor in relation to ratios

6

X(0 ) X(1)

7

and

R(0) X(1)

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o example

Let us consider a YNyn, 33 kV/11 kV transformer with a power rating of Sn = 24 MVA (see IEC 909-2 table 3 A) supplying a network with 240 mm² aluminium cables the longest outgoing feeder of which is 5 km. The neutral earth electrode resistance is 0.5 Ω. - transformer characteristics: Usc = 24.2 % RT = 0.046 XT X(0 )T XT

we can deduce

= 0.7

(

)

11 × 10 3 U2 XT = Usc × n = 0.242 × = 1.22 Ω Sn 24 × 10 6

RT = 0.056 Ω X(0 )T = 0.85 Ω Note:

the value of Usc is extremely high in relation to the transformers feeding a network with a limiting resistor earthed neutral. The transformer here is a United Kingdom transformer adapted to the solidly earthed neutral system.

The short-circuit voltage has been chosen high on purpose so as to minimise the short-circuit R(0) is minimised since X(1) = XT + XC , which current. Indeed, if Usc is high, the value X(1)

(

)

decreases the overvoltage factor (see fig. 5-3 and 5-4). - cable characteristics: RC =

ρ L 0.036 × 1000 = = 0.15 Ω / km S 240

XC = 0.1 Ω / km We assume that X(0 )C = 3 XC = 0.3 Ω / km . Note:

the value of X( 0 )C is highly variable (from 0.2 to 4 X(1) ) depending on what the cable is made of and the return via the earth (remote earth, screen or earthing conductor).

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For a solid fault ( R f = 0 ) at the transformer terminals: R(1) = RT = 0.056 Ω R(0) = 3 Re + RT = 3 × 0.5 + 0.056 = 1.56 Ω X(1) = XT = 1.22 Ω X(0 ) = X(0 )T = 0.85 Ω whence

R(1) = 0.05 X(1) ≅ 0 R(0) X(1) X(0 ) X(1)

= 1.28

= 0.70

Figure 5-3 shows that k is between 1.4 and 1.5. For a solid fault ( R f = 0 ) 5 km away from the transformer: R(1) = RT + RC = 0.056 + 0.15 × 5 = 0.81 Ω R(0) = 3 Re + RT + RC = 3 × 0.5 + 0.056 + 0.15 × 5 = 2.31 Ω X(1) = XT + XC = 1.22 + 0.1 × 5 = 1.72 Ω X(0 ) = X(0 )T + X(0 )C = 0.85 + 0.3 × 5 = 2.35 Ω whence

R(1) = 0.47 X(1) R(0) X(1) X( 0) X(1)

= 1.34

= 1.37

Figure 5-4 shows that k is between 1.2 and 1.3.

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n TN earthing system The current of an earth fault circulates in the protective conductor. The neutral earth electrode resistance is thus not used to determine the zero-sequence impedance (see fig. 5-5). ZT

ZC

ZT

ZC

ZT

ZC

V3 V2

V3

V1

V2

Z PE

Re

V1 , V2 , V3 ZT ZC Z PE VM Re

VM

VM

VM

: single-phase voltages : transformer impedance : cable impedance : protective conductor impedance : potential of exposed conductive parts (masses) in relation to earth : neutral earth electrode resistance

Figure 5-5: equivalent diagram of an earth fault in a TN earthing system

We are interested in the overvoltage of the healthy phases in relation to the exposed conductive part, which determines whether or not an insulation fault may occur on the other V − VM V − VM load: k M = 2 = 3 . Vn Vn

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For a transformer or a cable in low voltage, we can take the zero-sequence impedance to be approximately equal to the positive-sequence impedance: Z(0)T = ZT and Z(0)C = ZC . We thus have Z(0) = Z T + ZC + 3 Z PE Z(1) = ZT + ZC

whence

a=e

j 2π 3

kM = 1−

a 3 Z PE a Z PE for a solid fault ( R f = 0 ) = 1− 3 ( ZT + Z C + Z PE ) Z PE + ZT + Z C

: rotation operator of 120°

The overvoltage will be maximum when Z T is negligible compared with Z PE + ZC , which is the case for a long length cable. Thus

kM ≤ 1−

a Z PE Z PE + ZC

k M will be maximum when the protective conductor cross-sectional area is as small as possible, i.e. equal to half the phase conductor cross-sectional area; thus RPE = 2 RC . For an aluminium cable cross-sectional area smaller than 120 mm², the reactance can be neglected compared with the resistance, which thus gives us: 2 Z PE RPE ≅ = Z PE + ZC RPE + RC 3 whence

kM ≤ 1−

2 a 3

kM ≤ 1−

2 1 3  − + j 3 2 2 

since RPE = 2 RC

k M ≤ 1.45 We can show that for a cable with a large cross-sectional area (> 120 mm²), the overvoltage will be lower than in the case of a small cross-sectional area.

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n TT earthing system (see fig. 5-6) ZT

ZC

ZT

ZC

ZT

ZC

V3 V2 V1 If load 1

If

VN

Re R M1 RM 2 V M1

VM1

Re

load 2

RM1

R M2

: substation earth electrode resistance : load 1 and fault load earth electrode resistance : load 2 earth electrode resistance : load 1 and fault load phase-to-earth voltage

Figure 5-6: equivalent diagram of an earth fault in a TT earthing system

We want to know the overvoltage of the healthy phases in relation to the exposed conductive part, which determines whether or not an insulation fault may occur on the other load: V − VM V − VM kM = 2 = 3 Vn Vn In low voltage, the neutral and load earth electrode resistances are very high in relation to the transformer and cable impedance ( Z T and Z C are roughly several tens of mΩ). We can thus write that the fault current is: If =

V1 Re + RM1

and

( ZT + ZC ) I f

≅0

The exposed conductive part of load 1 is connected to phase 1 by the fault (zero impedance). The voltage of one healthy phase of this load in relation to the frame is V2 − V1 or V3 − V1

(since ( Z T + ZC ) I f ≅ 0 ) , whence k M = 3 = 1.73 .

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The exposed conductive part of load 2 is at the same potential as the remote earth. The voltage of one healthy phase of this load in relation to the exposed conductive part is therefore V2 − VNeutral or V3 − VNeutral : V2 − VNeutral = V2 − Re I f = V2 −

 a Re  Re Re V1 = V2 − a V2 = V2 1 −  Re + RM  Re + RM Re + RM 

a Re Re + RM

whence

kM = 1 −

for

RM = Re , k M = 1.32

for

RM > Re , k M < 1.32

The earth electrode resistance of a group of loads is in general higher than the substation earth electrode resistance. The overvoltage coefficient will thus be lower than 1.32 on load 2. The overvoltage factor is maximum in the TT earthing system for a load having an exposed conductive part connected to the same earth electrode as the fault load, we thus have kM = 3

n recapitulative table of maximum earth fault overvoltages in relation to the neutral earthing system

Medium and high voltage (1)

Low voltage (2)

solidly earthed neutral (HV or MV)

unearthed or impedance earthed neutral (MV)

TN system

TT system

IT system

< 1.73 * (generally 1.2 to 1.4)

1.73

1.45

1.73

1.73

(1) : phase-earth overvoltage (2) : phase-exposed-conductive-part overvoltage (*) : a network with a solidly earthed neutral is generally made up so as to limit overvoltages to values close to 1.2 to 1.4.

Table 5-1: maximum overvoltage factor in relation to neutral earthing system Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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n consequences on equipment selection The overvoltage factor and fault duration influence the choice of equipment insulation voltage level.

o solidly or limiting impedance earthed neutral in MV, or TT

and TN earthing system

in LV Rapid clearance of the fault, and thus a short overvoltage time, means that the switchgear phase-earth insulation level does not have to be higher than the nominal single-phase voltage.

o unearthed neutral in MV or IT earthing system in LV

Since the power supply does not have to be interrupted on occurrence of a first fault, the overvoltage is likely to occur for a long period of time (several hours). It is therefore advisable to choose switchgear with a phase-earth insulation level that is suitable for the nominal phaseto-phase voltage. Note:

some manufacturers give a phase-earth insulation withstand equal to the single-phase voltage, but stipulate that their switchgear can be implemented in an unearthed neutral network. There are also switchgear standards that specify an insulation level compatible with use in an unearthed neutral network.

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5.1.2.2.

Resonance and ferro-resonance

resonance The presence of inductive L , capacitive C and resistive R elements, connected, either in series or in parallel, causes spreading of current and voltage having values which may be dangerous for equipment.

series resonance Figure 5-7 shows a series R , L, C circuit at the terminals of which a voltage U is applied. I

R

L

C

U

Figure 5-7: series R , L, C circuit fed by a voltage U

The voltage U is the vectorial sum of the voltages at the terminals of each element: U

= U R + U L + UC = R I + j Lω I +

1 jCω

The vectorial diagram in figure 5-8 shows that for certain values of L and C , the voltages at the terminals of the inductance and capacitance may be higher than the network voltage U :

jL

I 1 jC

I

RI

U

Figure 5-8: vectorial diagram of a series R , L, C circuit fed by a voltage U

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The resonance phenomenon occurs when U L = − U C : j Lω I = −

1 j Cω

LC ω 2 = 1 We thus have U = R I ; the series inductance and capacitance behave like a short circuit. For given values of L and C , the angular frequency ω r such that LC ω 2r = 1 is said to be a resonant angular frequency. An overvoltage factor

f

is thus defined which is the ratio of the voltage U L (or U C ) to the

supply voltage U : f =

U L Lω r I = U RI

f =

Lω r 1 = R RC ω r

parallel resonance Figure 5-9 shows a parallel

R , L, C

circuit at the terminals of which a current source J is

applied.

IR

IL

R

L

IC C

J

U

Figure 5-9: parallel R , L, C circuit fed by a current source J

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The voltage U is common to the three elements. We have the following relation: 1  1 J = + + j Cω U  R j Lω  The resonance phenomenon occurs when I L = − IC : U = − jCωU j Lω L C ω2 =1

We thus have U = R J ; the inductance and capacitance behave like an open circuit. For given values of L and C , the angular frequency ω r such that LC ω 2r = 1 is said to be a resonant angular frequency.

An overvoltage factor is thus defined which is the ratio: - between the voltage that is produced at the terminals of the parallel R , L, C circuit when the resonance occurs - and the voltage that would be produced on occurrence of the resonance if the inductance (or capacitance) were the only circuit element f =

RJ Lω r J

f =

R = RCωr Lω r

The most current example of parallel resonance is the case of a network having harmonic currents (patterned by current sources) and reactive energy compensation capacitors (see § 8.1.5).

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example: resonance in a Petersen coil earthed HV/MV substation Figure 5-10 shows the diagram of a Petersen coil earthed HV/MV substation when an HV earth fault flows through the common earth electrode. HV

HV / MV transformer

If

C

Lc , Rc

Re1 Ve

If

Z MV

Re

: HV earth fault current

Lc , Rc : Petersen coil inductance and resistance Re , Re1 : earth electrode resistances C Ve Z MV

: MV cable phase-earth capacitance : rise in substation earth potential : sum of MV cable and transformer impedances

Figure 5-10: HV earth fault in an HV/MV substation with a Petersen coil earthed neutral

The symmetrical component method gives us the fault current value as (see § 4.2.2 of the Network protection guide): If =

where

3 Vn Z(1) + Z( 2 ) + Z( 0 )

Z(1) = ZT + Z l Z( 2 ) = Z T + Z l (0

Z( )T + Z 0)l +

e

Re1

ZT , Z( 0)T : HV transformer positive-sequence (or negative-sequence) and zero-sequence impedances Z l , Z(0 )l : HV line positive-sequence (or negative-sequence) and zero-sequence impedances

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In high voltage, the substation earth electrode value ( Re ) is very low compared with the transformer and line impedances. The fault current is thus independent of Re ; it is thus considered to be a source of current with a value of I f . The equivalent Thevenin’s diagram of the current source I f Re is shown in figure 5-11.

with an internal impedance of

Re equivalent If

Re

Ve

Re I f

Figure 5-11: equivalent Thevenin’s diagram of the current source I f with an internal impedance of Re

The equivalent MV network diagram is thus that shown in figure 5-12. Re

Ve

Rc

Lc

Z MV

Z MV

C

C

Z MV

Re I f C

Figure 5-12: equivalent MV network diagram on occurrence of an earth fault on the substation HV side

The transformer and cable impedances are negligible compared with the cable phase-earth 1 . capacitance: Z MV << Cω

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The simplified MV network diagram is thus that shown in figure 5-13. Re

Rc

Lc VL

Ve

Re I f

3C

VC

Figure 5-13: simplified diagram

Let VL be the voltage at the inductance terminals. We have

VL =

Lc ω

(

Re + Rc + j Lc ω − 3 C1ω

)

Ve

In the case of a Petersen coil earthed neutral, (resonance) tuning between the inductance and 1 and the MV cable capacitance is aimed at as far as possible. We thus have : Lc ω ≈ 3Cω VC ≈ VL whence VL = Lc ω Ve . Re + Rc

To minimise the rise in substation earth potential (Ve ), the resistance earth electrode must be as weak as possible (of the order of 0.5 Ω). We can thus neglect Re compared with Rc , which thus gives us: L ω VL = VC = c Ve = Q Ve Rc VC = Q Re I f

Q : coil quality factor VC : is equal to the MV cable phase-earth overvoltage in this case The coil quality factor must not therefore be too high in order to avoid the risk of a very high overvoltage. This is why, in some cases, a resistor must be connected in parallel with the coil, in order to reduce the quality factor.

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Numerical application: Let us take a 63/5.5 kV substation where: Vn =

5.5 = 3175 . kV 3

I f (63 kV ) = 3 kA Re = 0.5 Ω L ω Q= c =4 Rc The rise in potential is: Ve = Re × I f = 1 500 V . The phase-earth voltage in the cables is: VC = Q × Ve . VC = 1500 × 4 = 6 000 V VC = 1.89 Vn

The overvoltage in the cables is roughly twice the nominal phase-earth voltage. It can be dangerous if the substation earth electrode is of poor quality. Indeed, for Re = 3 Ω we will have VC = 11.3 Vn . It is thus essential to limit the value of Re .

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n ferro-resonance o parallel ferro-resonance (see fig. 5-14)

Let us take a circuit made up of a parallel-connected capacitance, a coil with a saturable iron core and a resistance. Let R be the resistance, C the capacitance and L the inductance which varies with the current flowing through the coil and the voltage at the circuit terminals. IT IL

IR R

IC

L

C

V

Figure 5-14: parallel ferro-resonance

The total current IT flowing through the circuit is then given by the relation (1): IT =

V + j (C ω V − I L ) R

(1)

We cannot express I L as a function of V , owing to the saturation. The rms values are given by the relation (2): IT2 =

V2 R

2

+ (C ω V − I L )

2

(2)

We can thus write relation (3) as follows: I T2 −

V2 R

2

= C ω V − IL

(3)

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This equation can be graphically resolved by plotting, as a function of V, the curves representing functions (see fig. 5-15): I = IT2 −

V2 R2

I = C ω V − IL

(a)

(b)

For any value of IT , the intersection of curves (a) and (b) gives the V solutions of equation (3); figure 5-15 shows the graphic resolution of this equation. Curve (a) is an ellipse having the equation: V2 R

2

+ I 2 = IT2

and having one half axis which is equal to IT and the other to R IT . An ellipse corresponds to each total current value IT . Curve I L (V ) presents a very steep slope when V increases owing to the saturation of the V . coil's iron core: I L (V ) = L (V ) ω

On saturation, L (V ) becomes very weak and the current then highly increases (see fig. 5-15). Curve I C = C ω V is a linear function of V (see fig. 5-15). Curve (b) shows the development of I C − I L = (C ω V − I L ) as a function of the voltage.

The OSA portion of curve (b) corresponds to a lead current in relation to the voltage owing to the preponderance of the capacitive current. On the other hand, the AB part corresponds to a lag current, since the inductive current is preponderant. The intersection of ellipse (a) and curve (b) can give: - an operating point Q if ellipse (a) is inside ellipse (a") passing through point A - three operating points M , N , P if ellipse (a) is between ellipses (a') and (a") - two points S , T if ellipse (a) is equal to ellipse (a') - a single point X if ellipse (a) is outside ellipse (a'). The ferro-resonant mechanism is described below.

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With the circuit being initially unused, the total current IT is zero, as well as the voltage V , and ellipse (a) is reduced to point O . If the current increases, the length of the axes of ellipse (a) increases and the voltage rises, the operating point M moves along branch OS of curve (b). When the total current exceeds the value I T' for which ellipse (a') cuts curve (b) at S , the operating point suddenly jumps from point M to point T located on branch AB of curve (b), it then moves along this branch. The voltage thus suddenly increases, going from VS to VT , and then it continues to increase if the current IT increases. If the total current now decreases, the operating point moves along branch AB and stays there, even if the current drops below the value I T' corresponding to ellipse (a'). When the current reaches the value IT , the operating point is P instead of M . It only returns to branch OS if the current drops below the value I T'' corresponding to ellipse (a") passing through point A . When this occurs, the operating point suddenly jumps from A to Q , and the voltage from V A to VQ . We can thus see that two stable operating conditions, for which the voltage at the circuit terminals takes very different values, for example V M and VP , can correspond to the same rms current value IT . Finally, if the initial operating conditions correspond to a weak voltage (branch OS ), with a resulting capacitive current, it is possible that, following a sudden change in operating conditions leading to a transient phenomenon (overcurrent or overvoltage), the resulting current becomes inductive and the voltage maintains a high value, even once the disturbance has disappeared. Ferro-resonance can be avoided if the resistance R is sufficiently weak for ellipse (a) to remain within zone OSA , even when there is a high overcurrent. IL

I

inductive operating conditions

IC

C V (b)

resonance

capacitive operating conditions

C V B (a''')

X

IT'''

S

'

IT IT IT''

Q

O

VQ

M

T N

(a'')

VM

IL

P

(a') (a)

A VS V N VA V P VT

V

Figure 5-15: parallel ferro-resonance - graphic resolution Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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o series ferro-resonance (see fig. 5-16)

Let us take a series circuit made up of a resistance, a coil with a saturable iron core and a capacitance. We have: I   V = R I + j  VL −   Cω 

(1)

We cannot express VL as a function of I , owing to the saturation. If we move to rms values, we can write:

or:

I   V 2 = R 2 I 2 + VL −   Cω 

2

I   V 2 − R2 I 2 = VL −   Cω 

2

(2)

(3)

I V 2 − R2 I 2 = Lω I − Cω

(4)

R VR

I

L

C

VL

VC

V Figure 5-16: series ferro-resonance

As for the parallel circuit, this equation can be graphically resolved as a function of I , by plotting curves (see fig. 5-17): v = V 2 − R2 I 2 and

v = VL −

I Cω

Curve VL ( I ) presents a very small slope when I increases owing to the saturation of the coil's iron core VL ( I ) = L ( I ) ω V .

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On saturation, L ( I ) becomes very weak and the voltage almost stops increasing when

I

rises. The network operating point is located at the intersection of curve (b) having the equation: v = VL −

I Cω

and ellipse (a) having the equation: v = V 2 − R2 I 2 There are three possible operating points: M , N , P . M and P are stable, N is unstable. A voltage disturbance can make the circuit move from point M to point P . This results in a high current and high overvoltages at the inductance and capacitance terminals. Ferroresonance can be avoided if the resistance R is sufficiently high for ellipse (a) to stay within zone OSA , even when there is a high overvoltage. V

VL

resonance

VC

V ''' V

(a''')

'

V

(a')

S N

V Q O

(a)

M

''

IQ

(b)

T P

A

(a'')

IM

X

IS

IN IA

IP

IT

IX

I

Figure 5-17: series ferro-resonance - graphic resolution

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o example of parallel ferro-resonance - unearthed neutral three-phase network (see fig. 5-18)

Let us consider a three-phase network with unearthed neutral having a capacitance C between each phase and the earth. Furthermore, a voltage transformer, with a similar magnetizing inductance to a saturable core reactor, is connected between each phase and earth. A parallel inductance-capacitance circuit thus appears between each phase and earth. Parallel ferro-resonance can then be sparked between the capacitance and voltage transformer of the same phase. This ferro-resonance may occur following a transient overcurrent or overvoltage caused by a switching operation and notably when the network is energized. Owing to the existing phase displacements between the voltages of the three network conductors, the overcurrents and switching overvoltages do not have the same magnitude in the three phases. Ferro-resonance can thus very easily occur on only two phases, phases 2 and 3 for example. The voltages of these two phases in relation to earth correspond to points located on portion AB of curve (b) (see fig. 5-15). The voltage of phase 1 corresponds to a point located on the OS part of this curve.

For phases 2 and 3, the capacitance-inductance assembly behaves like an inductance, and for phase 1, like a capacitance. If we plot the voltage vector diagram, we can see: - that the phase 1 voltage in relation to earth is weak - that the voltages in relation to earth of the other two phases are very high - that there is a very high potential difference between the neutral point and earth (see fig. 5-18).

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These overvoltages will cause a breakdown in equipment insulation if provisions to limit them are not taken. V3 Ph 3

V2 Ph 2

V1 Ph 1

N C

v1T

C L

C L

L

V1 vN v3T

v2T N V3

V2

Ferro-resonance occuring between two phases Figure 5-18: parallel ferro-resonance in an unearthed neutral network

Ÿ protection against the risks of parallel ferro-resonance A voltage transformer (VT ) charged by a resistor r behaves like a saturable (magnetizing) inductor in parallel with this resistor. Thus, in an unearthed network, if a charging resistor is connected to the secondary of the voltage transformers, the L-C parallel circuits, made up of these transformers and network cable capacitances, are transformed into R-L-C parallel circuits, such that if the resistors are correctly sized, the risk of ferro-resonance outlined previously can be avoided (ellipse (a) remains inside the zone 0SA - see fig. 5-15): - the resistors must be sufficiently weak to be efficient - they must not be too weak, so that the maintained.

VT

are not overcharged and their accuracy is

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In the case of VT (see fig. 5-19).

with a single secondary, a charging resistor is installed on each phase

A resistance value equal to 68 Ω is recommended for a secondary voltage of

VT

VT

100 V . 3

VT

r r r

measurements

Figure 5-19: protection against risks of ferro-resonance using resistors with single secondary VT

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In the case of VT with two secondaries, a resistor is installed in the open delta of one of the two (see fig. 5-20). It is recommended that a power above 50 W be dissipated in the resistor on occurrence of a phase-earth fault. 100 V , on occurrence of a solid earth fault, the voltage at the 3 resistor terminals is equal to 100 V; the resistance value is then determined: For a secondary voltage of

R≤

(100) 2 50

R ≤ 200 Ω

VT

VT

VT

measurements

r

Figure 5-20: protection against the risks of ferro-resonance via a resistor with two-secondary VT

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o example of series ferro-resonance (see fig. 5-21)

Figure 5-21 shows a solidly earthed network feeding a three-phase transformer having a deltaconnected primary. This can also apply to a star-connected transformer with an unearthed neutral. If, when the switch is closed, one of the poles remains accidentally open or closes late, for example the pole of phase 1, series ferro-resonance may occur in the circuit including: - the magnetizing inductance of transformer windings AC or BC - the capacitance of phase 1 in relation to earth.

Very high overvoltages can occur at the transformer terminals and between phase 1 and the earth. This type of ferro-resonance has frequently been encountered on HV networks with solidly earthed neutral. It may also occur when a switch is opened. The means of protecting against this type of ferro-resonance consists in inserting a resistor in the supply transformer neutral point earthing. This solution does not however provide total protection since ferro-resonance can, for example, occur in the circuit including the transformer AC winding and the capacitances of phases 1 and 3 in relation to earth. V3

switch

V2

Ph 3

V1

Ph 2

A L B

Ph 1

C

L

C

If

C

L

C

Figure 5-21: series ferro-resonance

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5.1.2.3.

Neutral conductor breakdown

Let us consider the diagram in figure 5-22 , where Z1 , Z 2 and Z 3 represent the equivalent impedances per phase of all the loads downstream of the neutral breakdown point. If the phases are perfectly balanced, the voltage system is not disturbed. In the event of load unbalance, the neutral point is displaced and the phase-neutral voltages move close to the phase-to-phase voltage for the least loaded phases, while for the loaded phases (weak impedance), they drop below the single-phase voltage. Z3 V3 Z2 V2 Z1 V1

N

neutral breakdown

Figure 5-22: equivalent diagram of an LV network during neutral breakdown

Using the superposition theorem, we can show that:  Z2 // Z3   Z1 // Z 3   Z1 // Z2  VN =   V1  V2 +   V3 +   Z1 + Z 2 // Z3   Z2 + Z1 // Z 3   Z3 + Z1 // Z2 

(1)

The voltage applied to the terminals of a single-phase load on phase 3, for example, will be: V3 N = V3 − V N If we know that V2 = a 2 V1

and

 1 3 V3 = a V1 ,  a = − + j  2 2  

then we can calculate V3 N , for example, for the following impedances: Z1 = R Z2 = 2 R Z 3 = 10 R (We have taken resistive loads to simplify the calculations.)

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By applying formula (1), we find:

 4 a2 + 9   V1 VN =   16 

we then have V3 N = V3 − VN = a V1 − VN

 − 15 + j 10 3  V3 N =   V1 16  

whence

V3 N = 1.43 Vn

Similarly, we can determine: V2 N = 114 . Vn and

V1N = 0.6 Vn

Vn : nominal single-phase voltage

We can see that once the most sensitive single-phase loads have broken down, there are successive breakdowns, following the development of the phenomenon which worsens the unbalance ( Z 3 increases after the breakdowns and consequently V3 N increases); this is an avalanche phenomenon. This risk thus underlines that it is preferable to well balance the loads on the three phases.

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5.1.3.

Switching overvoltages

When switching to energize or de-energize loads transient overvoltages occur on the network. These overvoltages are all the more dangerous if the current interrupted is inductive or capacitive. The magnitude, frequency and damping duration of these transient overvoltages depend on the given network characteristics and the mechanical and dielectric characteristics of the switching device. 5.1.3.1.

Interrupting principle

Interrupting an electric current using an ideal device involves the resistance of the device going from zero before interruption to an infinite value just after interruption. The interruption occurs the instant the current crosses zero. It is impossible to make such an ideal device, but with the interrupting techniques being based on the behaviour of the electric arc in different dielectric media we can come close to it. n circuit-breaker interruption The instant the current is interrupted, an electric arc is created between the terminals of the switching device. The conductive electric arc tends to be held by the ionizing phenomenon of the dielectric caused by the energy dissipated. Around current zero crossing, the dissipated energy decreases dropping below the thermal energy supplied to the medium, the arc cools down and its resistance increases. When the current crosses zero, the arc resistance becomes infinite and the arc is interrupted. Between the start and end of interruption, the voltage between the poles of the switching device goes from zero to the network voltage. This change gives rise to a high frequency transient phenomenon called the transient recovery voltage (see fig. 5-23).

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R

L I

V

VA

C

I

V

VA

t

L , R : inductance and resistance equivalent to the network upstream of the circuit-breaker C

: upstream network capacitance

Figure 5-23: transient recovery voltage during circuit-breaker interruption

n fuse interruption On occurrence of a short circuit, the value of the current flowing through the fuse is higher than its nominal fusing value. Interruption can thus occur at any instant and not necessarily the moment the current crosses zero. Figure 5-24 gives an example of a transient overvoltage which occurs on the network after a wire fuse has fused. Volts 1000

225

t ~ 1 ms

Figure 5-24: transient overvoltage on fusion of a wire fuse Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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5.1.3.2.

Load switching

n de-energizing loads o inductive load

Ÿ single-phase circuit Let us consider the equivalent single-phase diagram in figure 5-25 with an ideal circuit-breaker CB which has a zero arc resistance the instant the contacts separate and which carries out interruption when the current crosses zero. Before operation of the circuit-breaker, between points A and B, there is a voltage drop due to the load current flowing through Ls . At the instant of interruption, the voltage at B suddenly reaches the voltage at A and the capacitance Cs is charged through Ls . The energy exchanges between Cs and Ls make voltage oscillations at frequencies of 5 to 10 kHz occur. The voltage at C suddenly decreases to zero and the capacitance C p is then discharged through L . The energy exchanges between

and

Cp

create voltage oscillations at

L

frequencies going from 1 to 100 KHz. CB B ID

A Is

C

IL

Ls VA

Cs

L

Cp I0 Lp

Ls : network inductance upstream of the circuit-breaker Cs : network capacitance upstream of the circuit-breaker L : load inductance L p : stray inductance C p : network capacitance downstream of the circuit-breaker CB : circuit-breaker Figure 5-25: interruption in an inductive load network

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The phenomena observed are illustrated by curves in figure 5-26.

VA t

VB t

VC t Is t

ID t

IL t VD = VB − VC t t0 t0 t1

t1

: separation of contacts : zero current

Figure 5-26: interruption cycle of an ideal device

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Ÿ three-phase circuit When the three-phase circuit in figure 5-27 is interrupted, the first phase which sees the current crossing zero interrupts this current. There follows a transient current circulating in the two uninterrupted phases. Thus, if phase 1 interrupts the current first a transient voltage is obtained between points C1 , C2 and earth which is capable of reaching a value of 2 V$n for an ideal circuit-breaker. For an actual circuit-breaker, the overvoltage coefficient is higher than or equal to 2. V$n : peak value of the phase-neutral nominal voltage Note:

the current crosses zero on the following phase after 1/3 of a period (7 ms at 50 Hz), while the period of oscillations is roughly 1 ms.

A1

V1

Ls

B1 Cs

A2

V2

Ls

A3

V3

B2

Ls

L2

C2

N Cp

Lp

B3 Cs

Cp

Lp

Cs

L1

C1

L3

C3

Lp

Cp

Figure 5-27: equivalent diagram of a three-phase circuit during interruption

Ÿ restrike phenomenon The instant a circuit is interrupted, the voltage at the terminals of the circuit-breaker quickly increases (roughly from 0.1 to 0.5 kV/µs). If the circuit-breaker poles separate shortly before the current reaches zero (for an inductive circuit, this corresponds to the maximum voltage), regeneration of the dielectric medium may not be sufficient to withstand the stress-voltage. Indeed, in this case, the voltage is maximum and the poles are closer together. Renewed breakdown then occurs accompanied by overvoltages with a peak to peak magnitude of 2 V$n . This phenomenon is called restrike. Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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Ÿ multiple restrike If we consider the single-phase diagram in figure 5-25, we can see that in the case of restrike, the voltage at point C almost instantaneously reaches the voltage at point B . The capacitance C p is charged by a high frequency current (roughly 1 MHz) circulating in the L p , Cs , CB and C p circuit. This high frequency current very quickly crosses zero (1 µs). If the circuit-breaker manages to interrupt the current at that moment, the restrike phenomenon is repeated as the distance between the circuit-breaker contacts is still very small. Furthermore, the peak-to-peak magnitude of the oscillation is then equal to 4 V$n . The overvoltage increase makes the occurrence of a second breakdown highly probable. Indeed, the increase in dielectric withstand through the increase in the distance between the circuit-breaker contacts may be lower than the increase in overvoltage. This is why a multiple restrike phenomenon occurs with overvoltages of increasing magnitude (see fig. 5-28). In theory, such a phenomenon may generate overvoltages having a peak value equal to the dielectric withstand limit of the open device, without a definite interruption of the current being obtained. In practice, this case remains exceptional as it is enough for one of the restrikes to allow the power frequency current to be restored; a new current half wave then flows through the circuit-breaker. The circuit-breaker interrupts this half-wave the moment it crosses zero when the distance between the contacts is sufficient. Thus the types of circuit-breakers undergoing multiple restrike usually manage to interrupt the current without causing overvoltages of very high magnitude. VC t

Figure 5-28: voltage VC in case of interruption with multiple restrike

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Ÿ chopping current (weak inductive currents) When weak currents, notably lower than the nominal current of the circuit-breaker, are interrupted, the arc which occurs occupies a small volume. It consequently undergoes considerable cooling linked to the circuit-breaker's capacity to interrupt much higher currents. Owing to this fact, the arc becomes unstable and its voltage may present relatively large variations, while its absolute value remains lower than the network voltage (case of SF6 or vacuum). These voltage variations may generate high frequency oscillating currents, with a magnitude that may reach 10% of the current at 50 Hz, in the nearby capacitances ( Cs , L p , C p circuit in figure 5-25). Superposing these high frequency currents on the current at 50 Hz results in multiple crossings of the current through zero around zero of the fundamental wave (see fig. 5-29). The circuit-breaker interrupts the current the first time it crosses zero while the load current (only the current at 50Hz) is not zero. The value of this current represents what we call the chopping current Ichop .

(

)

current in the circuit-breaker

I chop "chopping" current extinction possible

50 Hz wave

Figure 5-29: superposition of a high frequency oscillating current on a power frequency current

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The current is then interrupted as in the case in figure 5-25 except for the peak-to-peak  1 magnitude of the oscillations, due to the presence of energy stored in L  L I a2  which is  2 1  added to that in the capacitance C p  C p V$n2  . 2  If V$ is half the peak-to-peak maximum value of the oscillation at point C, we can write: c max

1 1 1 C p V$c2max = C p V$n2 + L I a2 2 2 2 L 2 I a in single-phase. V$c max = V$n2 + Cp V$n : phase-neutral nominal voltage peak value For a three-phase circuit V$n

must be added in order to take into account the transient

operating conditions linked to the non-simultaneous interruption of the phases, whence: L 2 Ia V$c max = V$n + V$n2 + Cp This phenomenon is notably problematic in the case of an arc furnace transformer power supply. Indeed, the transformer is generally connected not very far away from the busbar. Thus the value of C p is very weak and therefore the value of V$c max high. We can determine V$c max by taking: L : transformer leakage inductance C p : capacitance of the cable linking the circuit-breaker to the transformer I a : transformer magnetizing current Schneider carried out an analysis for a single-phase arc furnace transformer where: Vn = we find

15000 V ; 3

L = 8.26 H ;

C p = 14.75 nF ;

Ia = 4.36 A

V$c max = 8.5 V$n

Installing an R , C reduced to 2 V$ .

circuit in parallel with the circuit-breaker allowed the overvoltage to be

n

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Ÿ virtual chopping current - simultaneous interruption of the three phases The transients generated by the first phase that creates overvoltages may cause, owing to the capacitive coupling between the phases, oscillating currents inside circuits L p , C p, Cs of the other phases. It is thus possible to obtain zero current in these phases, immediately (several hundreds of a microsecond) after interruption of the first phase. If the circuit-breaker interrupts such currents, a chopping current phenomenon is then created with very high chopping current and overvoltage values.

Ÿ chopping current and multiple restrike Current chopping and multiple restrike are frequently linked. Overvoltages caused by current chopping can themselves lead to restrike. They are almost systematic in the case of the virtual chopping current.

o capacitive loads (see fig. 5-30)

Interruption of capacitive circuits, such as a capacitor bank or off-load cable, raises less difficulties than the interruption of inductive circuits. Indeed, the capacitances remain charged at the peak value of the 50 Hz wave after extinction of the arc when the current reaches zero and the recurrence of voltage at the switchgear terminals is accompanied by a 50 Hz wave. Nevertheless, one half period after interruption, the device is subjected to a voltage equal to twice the 50 Hz peak voltage 2 V$n .

( )

If the speed and dielectric withstand of the device are not sufficient to withstand this stress, restrike may occur. It is followed by a voltage reversal at the terminals of the capacitances, raising them to a phase-neutral voltage equal to 3 V$n maximum (if damping is neglected).

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When the supply voltage reverses back, a half period later, the potential difference at the device terminals then reaches 4 V$n . Such an overvoltage can obviously cause renewed restrike between the device contacts, and the previously described oscillation mechanism is renewed with increased magnitude, leading to a new rise in the phase-neutral voltage of the capacitances 5 V$ .

( n)

The cumulative effect of multiple restrike is obviously highly dangerous for the network components as for the device itself. This rise in overvoltages can be avoided by choosing the appropriate equipment, i.e. which does not allow restrike.

VC

5V$n

20 ms

V

V$n

VC V$n

2 V$n t 4 V$n interruption

VC

3V$n

Figure 5-30: voltage rise on separation of a capacitor bank from the network by a slow operating device

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n energizing a load o inductive circuit

When a device closes, on an inductive circuit (no-load transformer, motor on starting), there is a moment when the dielectric withstand between its contacts drops below the applied voltage. A breakdown occurs causing sudden zero voltage at the device terminals. This is accompanied by oscillations with stray capacitances which cause high frequency currents to circulate in the circuit-breaker. Depending on the speed of the device, prestrikes may or may not occur up to complete closing of the poles. Multiple prestrike is accompanied by successive overvoltages which decrease until the device is completely closed. The phenomenon is highly complex and involves several parameters: - the characteristics of the switching device - the characteristic impedance of the connections - the natural frequencies of the load circuit which means that a mathematical simulation model is required to pre-determine the overvoltage values.

o capacitive circuit (capacitor bank)

When a capacitor bank is energized via a slow operating device, prestrike occurs between the contacts close to the wave peak of 50 Hz. A damped oscillation in the

system in figure 5-31 then occurs at a frequency above 50 Hz concentrated around the peak. In this case the maximum overvoltage is 2 V$n . It LC

corresponds to the maximum overvoltage admissible by the capacitors (see IEC 831-1 for LV and 871-1 for MV or HV). With a faster device, prestrike does not necessarily occur around the 50 Hz peak and consequently the overvoltage is smaller. When put out of service, the bank remains charged at a voltage going from 0 to the peak voltage of the network.

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If the bank is energized shortly afterwards, a breakdown due to the application of a voltage of opposite polarity may give rise to an overvoltage of 3 V$n .

L

CB

C

U

Figure 5-31: closing operation of a capacitive circuit

To ensure the safety of persons, the capacitor banks are fitted with a discharging resistor having a time constant allowing 75 V to be reached after 3 minutes in LV and 10 minutes in HV.

n Means of protecting loads The phenomena created by de-energizing (or energizing) loads, which we have studied, lead to transient overvoltages which may be dangerous for both loads and other network elements. Table 5-2 gives the level of overvoltages and their characteristics for each phenomenon studied. Occurrence of Number of phenomenon overvoltage peaks

Overvoltage value

dU/dt order of magnitude

Remark

Chopping current

at every interruption

2 to 4 V$n

0.1 kV/µs

favours restrike

Multiple restrike

interruption with 0 to 20 separation close to zero current

2 to 7 V$n

10 kV/µs

Prestrike

at every closing 1 to 50

2.5 V$n

10 kV/µs

1

V$n : phase-neutral voltage peak value Table 5-2: different types of overvoltage

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The loads affected by these phenomena are off-load transformers, neutral point coils (neutral reactance earthing) and motors during the starting period for inductive circuits as well as capacitor banks for capacitive circuits. Transformers undergo impulse wave dielectric tests; because of this, they are better built than motors to be able to withstand the transients caused by restrike (see IEC 76-3). The case of motors is different. At each start, they must withstand the transients caused by prestrike. Moreover, even if interruption during the starting period does not occur very often, it is nevertheless a possibility and they are then subjected to multiple restrike. Motors are thus especially sensitive to multiple prestrike, because of its high rate of occurrence, as well as to multiple restrike, due to the magnitude of the overvoltages produced. These overvoltages cause deterioration of the insulation of the first turns. In order to limit overvoltages, Zn0 type surge arresters can be connected in parallel with the load. But the best method consists in using switching devices suitable for the type of application. Table 5-3 gives the behaviour of medium voltage switchgear with respect to the phenomena relating to the switching overvoltages studied. Switchgear

Multiple prestrike on closing

Current chopping

Multiple restrike

no

weak

no

No problem. Below 300 kW, use a rotating arc SF6 circuit-breaker.

no

no

no

No problem.

Vacuum circuit-breaker

yes

yes

yes

Use surge arresters

Vacuum contactor

yes

weak

yes

Use surge arresters

Magnetic blast circuitbreaker and contactor

no

no

no

No problem.

Minimum oil circuit-breaker

no

yes

yes

Use surge arresters

Puffer-type SF6 circuitbreaker Rotating arc SF6 circuit-

Overall behaviour

breaker and contactor

Table 5-3: behaviour of medium voltage switchgear

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5.1.3.3.

Circuit-breaker clearance of phase-earth faults

Let us consider the three-phase network shown in figure 5-32 in which phase 1 is affected by an earth fault. In this case, the network is equivalent to the diagram in figure 5-33 which corresponds to the case examined in paragraph 5.1.2.1. At the start of contact separation, the arc voltage is weak and remains constant. On the other hand, just before interruption, this voltage, called the extinction voltage, increases to a more or less high value which may exceed V$n . This voltage depends on the type of circuit-breaker (air, oil, SF6 , vacuum) as well as the arc extinction technique (cooling, lengthening, rotating arc).

When the current crosses zero, the arc is extinguished and the recovery voltage magnitude will depend on the extinction voltage as follows: - for the case of neutral earthing via resistance (the fault current is in phase in relation to the voltage), the extinction voltage limits the magnitude of recovery voltage oscillations - for the case of neutral earthing via reactance (the fault current is phase shifted by

π in 2

relation to the voltage), the extinction voltage increases the magnitude of oscillations.

After interruption, restrike may take place if re-generation of the dielectric medium is not fast enough in relation to the rise in recovery voltage. In this case, the magnitude of oscillations may reach double the size of the first recovery voltage. If we neglect the transformer and line impedances, the voltage at the terminals of the neutral earthing impedance (VN ) is equal to the difference between the supply voltage and the voltage at the circuit-breaker terminals. The voltage V N is vectorially added to the voltage of the healthy phases and may lead to the latter reaching higher overvoltages than the overvoltages observed on the fault phase. The curves in figure 5-34 give the overvoltage levels recorded on occurrence of an earth fault in relation to the network characteristics and the earthing impedance.

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We can see that reactance earthing of the neutral (case with restrike) clearly increases the magnitude of the overvoltages. Resistance earthing is thus preferable. In the latter case, we see that the overvoltages do not exceed 240 % when the ratio of the current in the earthing resistor to the network capacitive current is equal to 2 (see fig. 5-34). In networks with resistance earthing, the following relation should therefore always be respected if possible: I rN > 2 I C I rN : current in the neutral earthing resistor during the fault I C : currents in the network phase-earth capacitances (see § 4.3 of Protection guide) V3

CB Ph 3

V2 Ph 2 V1 Ph 1 ZN or r N

C

ZN

: neutral earthing impedance (or

C If

: phase-earth capacitance : fault current

CB V1 , V2 , V3

: circuit-breaker

C

C

If

rN )

: single-phase voltages

Figure 5-32: phase-earth fault clearance CB

Xnet

~ ZN

C

or r N

If IC I rN

Xnet

: network reactance

C : fault phase earth capacitance Z N or rN : neutral earthing impedance (or resistance rN ) If : fault current Figure 5-33: fault circuit on occurrence of a phase-earth fault

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High resistance earthing with restrike in the fault or circuit-breaker, case of industrial networks for which IrN < 20 to 30 A (see Protection guide - § 10.1.1.). The overvoltage depends on the ratio

I rn IC

Limiting resistor earthing with restrike in the fault or circuit-breaker, case of public distribution networks for which I rN is equal to several hundred to 1 000 A. The overvoltage

rN Xd

depends on the ratio

transient voltage as % of the nominal single-phase voltage peak value

transient voltage as % of the nominal single-phase voltage peak value %

% 250

healthy phases

460 200 300 260 240 200

healthy phases

100

neutral fault phase 0.5

1

1.5

2

2.5

I rN IC

I rN = Vn / rN : current in the neutral resistor during the fault

I C = 3 Cω Vn : vectorial sum of current in the phase-

150

neutral

100

fault phase

50

3 4 5 6

8 10

20

30

50

70

90

rN X (1)

rN : neutral point earthing resistance X(1) : network positive-sequence reactance

earth capacitances

If I rN ≥ 2 I C , the overvoltage does not exceed 240 % Reactance earthing, case of public distribution networks for which I XN is equal to 1 000 to several thousand amps Case without restrike in the circuit-breaker

Case with restrike in the circuit-breaker

transient voltage as % of the nominal single-phase voltage peak value

transient voltage as % of the nominal single-phase voltage peak value %

%

500

400

400

B C

300

B

200

A

C

theoretical limits without damping

0

2

4

6

8

10 12 14 16

X(1) : network positive-sequence reactance X N : neutral point earthing reactance A : earth fault phase

theoretical limits without damping

A 200

N

N

100

300

100

XN X (1)

0

2

4

6

8

XN X (1)

10 12 14 16

B, C : healthy phases N : voltage at reactance terminals

Figure 5-34: transient overvoltages depending on the type of neutral earthing during a phase-earth fault Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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Reactance earthing: Voltage at circuit-breaker terminals

Resistance earthing: Voltage at circuit-breaker terminals

(network reactance) Voltage at the terminals of the reactance

(network reactance) Voltage at the terminals of the resistor

: arc extinction voltage

Figure 5-35: transient voltage on circuit-breaker opening during a permanent phase-earth fault

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5.1.4.

Atmospheric overvoltages

n general The earth and the electrosphere, the conductive area in the atmosphere (about 50 to 100 km thick), constitute a natural spherical capacitor which is charged by ionization, thus producing an electric field directed towards the ground of roughly several hundred volts/metre Since air is not very conductive, there is thus an associated permanent conduction current, of roughly 1 500 A for the entire earth's globe. Electrical balance is ensured during discharges by rain and strokes of lightning. The formation of storm clouds, masses of water in the form of aerosols, is accompanied by charge separation electrostatic phenomena: the positively charged light particles are driven by the rising air currents, and the negatively charged heavy particles fall because of their weight. At the base of the cloud, there may also be islets of positive charges where heavy rains are located. On an overall macroscopic scale, a dipole is created. When the breakdown withstand limit gradient is reached, a discharge is produced inside the cloud or between clouds or between the cloud and the ground. In the latter case, it is referred to as lightning. The cloud-ground electric field can reach 15 to 20 kV/metre on flat ground. But the presence of obstacles deforms and locally increases this field by a factor of 10 to 100 or even 1 000 depending on the form of the obstacles (also called the "peak effect"). The atmospheric air ionizing threshold is thus reached, i.e. roughly 30 kV/cm, and corona effect discharges are produced. When these discharges are located on fairly high objects (tower, chimney, pylon) they may divert lightning to this objects.

o classification and characteristics of strokes of lightning

Strokes of lightning are classed according to the origin of the discharge (or leader) and their polarity. Depending on the leader origin, the stroke of lightning may be: - either descending from the clouds to the ground in the case of fairly flat land - or ascending from the ground to the clouds in the case of mountainous land. Depending on the polarity the following distinctions between lightning strokes are made: - negative when the negative part of the cloud is discharged, which represents 80 % of cases in temperate countries - positive when the positive part of the cloud is discharged.

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Ÿ form and magnitude of the lightning wave The physical phenomenon of lightning corresponds to a source of impulse current the actual form of which is highly variable: it consists of a front rising up to the maximum magnitude of several miscroseconds to 20 µs followed by a decreasing tail of several tens of µs (see figure 5-36).

Figure 5-36: oscillogram of a lightning current

The magnitude of strokes of lightning varies according to a log-normal distribution law. We can thus determine the probability of a given magnitude being exceeded (see figure 5-37). We can see, for example, that for the average curve (IEEE), the probability of exceeding a magnitude of 100 kA is 5 %. This means that 95 % of lightning strokes have a magnitude less than 100 kA.

Figure 5-37: probability of exceeding positive and negative lightning stroke magnitudes, according to IEEE (experimental statistic)

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Similarly, the steepness of the wave front varies according to a log-normal distribution law. Let us determine the probability of exceeding a given front steepness (see fig. 5-38). We can see that the probability of exceeding a front steepness of 50 kA/µs of a negative stroke of lightning is 20 %.

Figure 5-38: probability of exceeding the front steepnesses of positive and negative lightning currents according to IEEE (experimental statistic)

Ÿ standard wave form The lightning impulse wave form given by IEC 71-1 is a 1.2/50 µs wave (see fig. 5-39): - rise time to the maximum value of 1.2 µs - time to half-value of 50 µs.

Figure 5-39: standard lightning impulse voltage wave form (IEC 71-1)

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Ÿ lightning density On a world-wide scale, 63 billion discharges are recorded on average each year which corresponds to 100 discharges per second. In France, this figure varies from 1.5 to 2 million lightning strokes per year. We then define the lightning density

as being the number of days per year on which

thunder has been heard in a place. In France, the average value of

is 20 with a variation range going from 10 in channel

coastal regions up to over 30 in mountainous regions. The value of

may be much higher and reach 180 in tropical Africa or Indonesia.

Ÿ lightning strike density The lightning strike density represents the number of lightning strikes per km2 per year, whatever their current value levels. In France,

varies between 2 and 6 lightning strikes/km2/year depending on the region.

o lightning impact mechanism

The lightning impact mechanism begins with a leader from a cloud which approaches the ground at a low speed. When the electric field is sufficient, sudden conduction is established giving rise to the lightning discharge. An experimental practical approach has enabled the relation linking the current of the lightning strike to the distance between the starting point (leader position) and discharge point (point of impact connected to the earth) to be found: or

according to the authors.

: striking distance, in m : lightning current, in kA

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By applying an electro-geometrical model to a vertical rod with a height we can show that there are two distinct zones:

(see fig. 5-40-a),

- zone 1 :

this is located between the ground and the parabola which is the locus of the equidistant points of and the ground; the instant the flash occurs, any leader located in this zone will touch the ground since it is nearer to this than to

- zone 2 :

this is located above the parabola; the instant the flash occurs, any leader located in this zone will be picked up by point on the vertical rod as soon as the distance between and the leader is less than the striking distance .

Figure 5-40-a: diagram of different protection zones offered by a vertical rod

For a lightning current with a value of , and thus a given striking distance, the distance x between the point of impact on the ground and the point where the rod is fixed to the ground (called the rod pick-up radius) is: if if The rod pick-up radius

is thus all the greater the more intensive the lightning stroke.

For very weak currents, the pick-up radius becomes less than the height of the rod which is then able to pick up the current along its length. This has been experimentally proved.

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Ÿ application to equipment protection using a lightning conductor The lightning conductor diverts lightning to itself in order to protect equipment. Its principle is based on the striking distance; tapered rods are placed at the top of equipment to be protected, they are connected to the earth by the most direct path (the lightning conductors surrounding the structure to be protected and interconnected with the earthing network). The electrogeometric model allows the zone to be protected to be determined using the fictive sphere method. The point of impact of the lightning is determined by the object on the ground the closest to the leader starting distance d . Everything happens as if the leader was surrounded by a fictive sphere with a radius d moving with it. For good protection, the fictive sphere rolling on the ground reaches the lightning conductor without touching the objects to be protected (see fig. 5-40-b). Protection against direct lightning strikes is approximately good in a cone the top of which is the top of the lightning conductor and the half-angle at the top is 45 °.

leader

d = critical striking distance

protected zone (cone)

fictive sphere

lightning conductor 45°

Figure 5-40-b: determining the zone protected by a lightning conductor using the "fictive" sphere method Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction , either partly of this document are not allowed without our written consent.

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n direct lightning strike (on phase conductors) When lightning strikes the phase conductor of a line, the current i (t ) is shared out in equal quantities on either side of the point of impact and is spread along the conductors. These have a wave impedance Z the value of which is between 300 and 500 Ω. This impedance is that seen by the wave front, is independent of the length of the line and of a different type from the impedance at 50 Hz. This results in a voltage wave of: U (t ) = Z .

i (t ) 2

which spreads along the line (see fig. 5-41).

U

i U =Z

i 2

i

t

Figure 5-41: lightning strike on a phase conductor

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Depending on the magnitude of the lightning current, two cases may occur:

o full impulse propagation

I   If the maximum voltage U max = Z max  is below the sparkover voltage U a of the insulator  2  string, the entire (full) wave spreads along the line.

o chopped impulse propagation

In the case where U max ≥ U a , as a first approximation, insulator sparkover occurs at the value of U a , and a phase-earth fault occurs at 50 Hz due to the arc being maintained. The lightning that is propagated is thus broken at the maximum value corresponding to U a . The lightning current causing this flashover, for a given line, is called the critical current

IC

such that: U IC = 2 a Z

For lines, the order of magnitude of IC is: - 5.5 kA at 225 kV, which corresponds to a probability of exceeding the magnitude according to the IEEE method of 95 % (see figure 5-37) - 8.5 kA at 400 kV, which corresponds to a probability of exceeding the magnitude according to the IEEE method of 92 % (see figure 5-37).

In medium voltage, flashover is systematic in the case of a stroke of lightning occurring due to the small distances in the air of the insulator string. This flashover of the insulator gives rise to a phase-earth fault current, called a follow current, which is held at the power frequency of 50 Hz until it is cleared by the protections.

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n indirect lightning strikes (lightning protection rope or pylons) When lightning strikes the line protection rope, part of the current flows through the pylon since the protection rope is connected to it (see fig. 5-42). This results in a potential rise at the top of the pylon the value of which depends on the self inductance L of the pylon and the resistance R of the earth electrode:  di ( t )  U (t ) = k  R i (t ) + L  dt   k

: ratio of the current shunted into the pylon by the incident current

lightning strike i k.i

U

protection rope

L

k .i

R

di   U = k R × I + L  dt   Figure 5-42: lightning strike on a protection rope

The voltage U may reach the impulse sparkover voltage of the insulators and cause a breakdown. This is "back-flashover". Part of the current is then propagated along the affected phase(s) towards the users. This current is in general greater than that of a direct lightning strike.

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In extra high voltage (> 220 kV), back-flashover is unlikely (the flashover level of the insulators is high), which is why it is useful to install protection ropes thus limiting the number of service interruptions. But below 90 kV back-flashover occurs even if the value of the earth electrode resistance is low (< 15 Ω); the usefulness of protection ropes is thus limited (more frequent service interruptions).

o induced impulse

A stroke of lightning that falls anywhere on the ground behaves like an electromagnetic field radiation source. The steeper the rising front of the lightning current the greater the radiation. For front steepnesses of 50 to 100 kA/µs, the effects of this field will be felt several hundreds of metres, if not kilometres, away. The magnetic field H at a point located at a distance of r current I flows, is given in the relation: H=

from a circuit through which a

I 2π r

This field creates induced voltages in the neighbouring circuits which are able to reach dangerous values both for equipment and persons.

Ÿ case of a loop Let us consider the loop formed by the supply cable and the telecommunication link in figure 5-43, with a surface S and located 100 m from the lightning impact which has a current rising front steepness of 80 kA/µs.

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The induced voltage is given in the relation: e=− µ 0 = 4 π × 10 −7

dφ dB dH = −S = − µ0 S dt dt dt

: magnetic permeability of the vacuum

now

103 1 dI 1 dH = = × 80 × −6 = 127 × 106 A / m / s dt 2 π r dt 2 π × 100 10

whence

e = 4 π 10 −7 × 120 × 127 × 106 = 19 kV

A phase-earth overvoltage of 19 kV thus occurs on the loop. This has a very short duration ( ≈ 1µs ) but can cause insulation breakdown. To avoid this risk, the surfaces of the loops must be reduced.

lightning impulse front steepness = 80 kA/µs

telecommunication link

computer

circuit loop surface = 120 m² supply cable

magnetic field 100 m printer

phase-earth insulation subjected to 19 kV ( 1 µs)

earth

Figure 5-43: circuit loop

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n impulse wave transference in a transformer (see IEC 71-2 - appendix A) In lightning impulse conditions, the transformer behaves like a capacitive divider with a ratio of s ≤ 0.4 . It is equivalent to a capacitance Ct (see figure 5-44-a).

equivalent

lightning wave

Ct U0

U1

U0

sU1

sU1

U1 : impulse voltage on the high voltage terminal U 0 : no-load voltage transferred Figure 5-44-a: impulse wave transference in a transformer

U 0 represents the no-load overvoltage, i.e. when the secondary outgoing terminals are not connected to any cables or lines. This overvoltage is generally not acceptable by the transformer. In reality, the transformer is connected to a network with a capacitiance Cs . This plays the role of a voltage divider with the transformer capacitance Ct (see fig. 5-44-b). Ct U 0 = sU1

Cs

U2

U 2 : voltage transferred to the secondary with a network Figure 5-44-b: transformer with its equivalent network

The voltage transferred to the secondary is thus:

U2 =

Ct s U1 Ct + Cs

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The values of Ct are generally between 1 and 10 nF. The capacitance of a cable is close to 0.4 nF/m. Thus, several tens of metres of cable will greatly reduce the overvoltage transferred to the secondary. In general, the network is sufficiently widespread for the overvoltage transferred not to raise any difficulties. However, in the case of a short cable, e.g. between a specific transformer and a load (arc furnace, etc.), the overvoltage transferred may be unacceptable for the equipment on the low voltage side. To reduce the magnitude of the impulse transferred, it is possible to: - use a surge arrester with a lower sparkover voltage on the high voltage side - install a surge arrester on the low voltage side between each phase and earth - increase the capacitance between each phase and earth on the low voltage side. 5.1.5.

Propagation of overvoltages

Overhead lines and cables constitute a propagation media for any overvoltage wave likely to occur on a network. For high frequencies (case of switching and lightning overvoltages), the line is characterised by its so-called "characteristic" or "wave" impedance: Zc ≈ L C

L C

: line inductance : line capacitance

We can see that this impedance is independent of the length of the line. The speed of the wave propagation on an overhead line is close to the speed of light: c = 3 × 108 m/s for cables, it is equal to v =

εr

(300 m/µs) c

εr

: relative permittivity of the cable insulating material

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The value of v is close to 150 m/µs. This gives us an idea of the way a lightning wave spreads along a conductor. Figure 5-45 shows how a lightning wave spreads along an overhead line in relation to time and space.

development in time V

front : 200 kV / µs

400 kV 2 µs

t (to x constant)

spread in space

V

front : 0.66 kV / m

400 kV 600 m = 300 x 2 µs

x (to t constant)

Figure 5-45: diagram showing how a lightning wave spreads along an overhead line in relation to time and space

Let us closely examine the phenomenon that is produced at a point M , where a change of impedance exists, separating two circuits with characterstic impedances of Z1 and Z 2 (see fig. 5-46).

v1 i1

Z1

v 1' ' i1

v2 i2

M

Z2

Z1 , Z 2 : upstream and downstream characteristic impedances v1 , i1 : incident wave upstream of M v2 , i2 : wave transferred downstream of M v1' , i1'

: wave reflected upstream of

M

Figure 5-46: propagation of a wave at a change of impedance point M Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction , either partly of this document are not allowed without our written consent.

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Upstream of M , we have:

v1 = Z1 i1 and v1' = − Z1 i1'

(1)

immediately downstream of M : v 2 = Z 2 i2

(2)

at point M : v2 = v1 + v1'

and i2 = i1 + i1'

(3)

We can thus deduce: v2 = v1 + v1' = v1 − Z1 i1' = v1 − Z1 ( i2 − i1 ) whence

v2 =

Z2 × 2 v1 Z 2 + Z1

In particular: - for a line short-circuited to earth, Z 2 = 0 ; we can deduce from this that v2 = 0 and v1' = − v1 - for a conductor without a change of impedance, Z 2 = Z1 ; we can deduce from this that v2 = v1 and v1' = 0 - for an open line, Z 2 = ∞ ; we can deduce from this that v2 = 2 v1 and v1' = v1 .

To conclude, at the point of change of impedance, the maximum voltage value may reach double the incident wave. This is the case of a line feeding a transformer as its impedance in relation to the lightning wave is very high in relation to the characteristic impedance of the line.

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5.2. 5.2.1.

Overvoltage protection devices Principle of protection

The protection of installations and persons against overvoltages is greatly improved when disturbances flow to earth, and this is done as close as possible to the sources of disturbance. This requires low impedance earth electrodes to be implemented. Thus, three overvoltage protection levels can be distinguished: n 1st protection level The objective is to avoid a direct impact on structures by catching the lightning and directing it towards designated flow points, via: - lightning conductors, whose principle is based on the striking distance; a rod placed at the top of a structure to be protected captures the lightning and evacuates it through the earthing network (see fig. 5-40-b) - meshed or Faraday cages - lightning protection ropes (see fig. 5-42). n 2nd protection level Its aim is to ensure that the basic impulse level (BIL) of the substation components has not been exceeded. In HV, this type of protection is established using elements ensuring that the lightning wave flows to earth, such as: - spark-gaps - HV surge arresters. n 3rd protection level Used in LV as an extra protection for sensitive equipment (computers, telecommunication devices, etc.). It uses: - series filters - overvoltage limiters - LV surge arresters.

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5.2.2.

Spark-gaps

n operation The spark-gap is a simple device made up of two electrodes, the first connected to the conductor to be protected and the second connected to earth. At the place where it is installed in the network, the spark-gap constitutes a weak point where overvoltages can flow to earth and thus protects the equipment. The sparkover voltage of the spark-gap is set by adjusting the distance in the air between the electrodes so as to obtain a margin between the impulse withstand of the equipment to be protected and the impulse sparkover voltage of the spark-gap (see fig. 5-47). For example, B = 40 mm on French public EDF 20 kV networks.

bird proof rod earth electrode

phase electrode

45°

45°

electrode holder

B

rigid anchoring chain device for adjusting B and locking the electrode

Figure 5-47: MV spark-gap with birdproof rod

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n advantages The main advantages of spark-gaps: - their low price - their simplicity - the possibility of setting the sparkover voltage.

n drawbacks - The sparkover characteristics of the spark-gap are highly variable (up to 40 %) depending on the atmospheric conditions (temperature, humidity, pressure) which modify the ionization of the dielectric medium (air) between the electrodes. - the sparkover level depends on the overvoltage. - spark-gap sparkover causes a power frequency phase-to-earth short circuit owing to the arc being maintained. The short circuit lasts until it is cleared by the switching devices (this short circuit is called a follow current). This means that it is necessary to install shunt circuitbreakers or rapid reclosing system on the circuit-breaker located upstream. Because of this, the spark-gaps are unsuitable for the protection of an installation against switching overvoltages. - the sparkover caused by a steep front overvoltage is not instantaneous. Due to this delay, the voltage actually reached in the network is higher than the chosen protection level. To take this phenomenon into account, it is necessary to study the voltage-time curves of the spark-gap. - sparkover causes the appearance of a steep front broken wave which could damage the windings of the transformers or motors located nearby. Although still used in certain public networks, spark-gaps are currently being replaced by surge arresters.

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5.2.3.

Surge arresters

To overcome the drawbacks of spark-gaps, different models of surge arresters have been designed with the aim of ensuring better protection of installations and good continuity of service. Non-linear resistor type gapped surge arresters are especially found in HV and MV installations which have been in operation for several years. The current tendency is to use zinc oxide surge arresters which provide better performance.

n definitions Surge arrester discharge current The surge or impulse current which flows through the arrester after a sparkover of the series gaps. Surge arrester follow current The current from the connected power source which flows through an arrester following the passage of discharge current. Surge arrester residual voltage The voltage that appears between the terminals of an arrester during the passage of discharge current. 5.2.3.1.

Non-linear resistor type gapped surge arresters (see IEC 99-1)

n operating principle In this type of surge arrester, a variable resistor (varistor), which limits the current after the passage of the impulse wave, is associated with a spark gap. After evacuation of the impulse wave to earth, the surge arrester is only subjected to the network voltage and the follow current is limited by the varistor. The arc is systematically extinguished after the 50 Hz wave of the single-phase-to-earth fault current has reached zero. Owing to the variation of the resistance, the residual voltage is maintained close to the sparkover level. Indeed, this resistance decreases with the increase in current. Various techniques have been used to make non-linear resistor type gapped arresters. The most conventional method uses a silicon carbide (SiC) resistor. Some surge arresters also have voltage grading systems (resistive or capacitive dividers) and arc blowing systems (magnets or blow-out coils).

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n characteristics

Variable resistor type surge arresters are characterised by: - the rated voltage, which is the maximum specified value of the power frequency rms voltage permitted between its terminals for which the surge arrester is designed to function correctly. This voltage can be continuously applied to the surge arrester without this modifying its operating characteristics. - the sparkover voltages for the different wave forms (power frequency, switching impulse, lightning impulse, etc.). - the impulse current evacuation capacity. 5.2.3.2.

Zinc oxide ( ZnO ) surge arresters

n operating principle Figure 5-48 shows that, unlike the non-linear resistor type gapped surge arrester, the zinc oxide surge arrester is only made up of a highly non-linear variable resistor. The resistance goes from 1.5 MΩ at the duty voltage (which corresponds to a leakage current below 10 mA) to 15 Ω during discharge. Following the passage of the discharge current, the voltage at the terminals of the surge arrester become equal to the network voltage. The current which flows through the surge arrester is very weak and is stabilised around the value of the earth leakage current. Because of the high non-linearity of the ZnO surge arrester a high current variation causes a low voltage variation (see fig. 5-49). For example, when the current is multiplied by 107, the voltage is only multiplied by 1.8.

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connecting spindle flange (aluminium alloy)

elastic stirrup rivet

exhaust pipe and overpressure device in the upper and lower flanges

Zn O blocks washer

fault indication plate spacer thermal shield exhaust pipe porcelain enclosure compression spring flange

rubber seal

prestressed tightness device

ring clamping device

overpressure device

Figure 5-48: example of the structure of a ZnO surge arrester in a porcelain enclosure for 20 kV networks peak kV U

600 500 400

Zn O

300 200 linear SiC 100

.001

.01

.1

1

10

100 1000 10000

I

SiC ZnO

: non-linear resistor type gapped surge arrester made up of a silicon carbide resistor : zinc oxide surge arrester

linear

:

U curve proportional to I

Figure 5-49: characteristics of two surge arresters having the same 550 kV/10 kA protection level Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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n characteristics ZnO surge arresters are characterised by: - the steady-state voltage which is the permitted specified value of the power frequency rms voltage that can be continuously applied between the terminals of the surge arrester - the rated voltage which is the maximum power frequency rms voltage permitted between its terminals for which the surge arrester is designed to operate correctly in the temporary overvoltage conditions defined in the operating tests (a power frequency overvoltage of 10 seconds is applied to the surge arrester - see IEC 99-4) - the protection level defined at random as being the residual voltage of the surge arrester when it is subjected to a given current impulse (5,10 or 20 kA according to the class), with a wave form of 8/20 µs - steep front current impulse (1 µs), lightning impulse (8/20 µs), long duration impulse, and switching impulse withstand - nominal discharge current.

Table 5-4 gives an example of the characteristics of a phase-to-earth ZnO surge arrester for a 20 kV public distribution network (with tripping on occurrence of the first fault).

Maximum steady-state voltage (phase-earth)

12.7 kV

Rated voltage Residual voltage for nominal discharge current

24 kV < 75 kV

Nominal discharge current (8/20 µs wave)

5 kA

Impulse current withstand (4/10 µs wave)

65 kA

Table 5-4: example of the characteristics of a ZnO surge arrester for a 20 kV network

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n choice of zinc oxide surge arresters in HV The general method for choosing a zinc oxide surge arrester in HV consists in determining its characteristic parameters using the network data, at the place where it will be installed. The parameters characterising the surge arrester are: -

U C , steady-state voltage

-

U r , rated voltage

-

I nd , nominal discharge current

- discharge class and energy capacity - mechanical characteristics.

The data relative to the network are: -

Um , highest phase-to-phase voltage applied to equipment

-

TOV temporary overvoltages (appearing on occurrence of an earth fault or load shedding on the public distribution network).

The choice of the surge arrester involves making a compromise between the equipment protection levels and the energy capacity of the surge arrester. The protection level must be as low as possible for the equipment withstand. This involves the lowest voltage rating possible and thus greater difficulty withstanding temporary overvoltages.

o determining U C and U r

Ÿ simplified method using equipment characteristics The voltages U C equipment Um : UC ≥

and U r

may be directly determined using the highest voltage for the

Um 3

Ur = 1.25 × UC

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Ÿ more accurate method using temporary overvoltages The simplified method has a drawback as it does not take into account the real requirements Um of the network which are generally lower than . 3 The temporary overvoltages likely to occur in a network are of two types: - overvoltages due to a phase-earth fault the clearance time of which depends on the protection system (see table 5.1 - the earth overvoltage factor is equal to 1.73 for unearthed or impedance earthed networks) - overvoltage due to load-shedding on the public distribution network, of the order of 15 % but able to reach 35 % in some networks. The temporary overvoltage value to be taken into account is the product of the earth fault overvoltage and load shedding factors. - specific case If one of the temporary overvoltages lasts over 2 hours, it is considered to be a steady-state condition for the surge arrester and thus U C is chosen to be equal to this overvoltage and Ur = 1.25 × UC - general case A surge arrester's capacity to withstand temporary overvoltages is given in relation to an equivalent voltage lasting 10 seconds (U10s ) expressed in the following equation: T U10s = TOV    10 

η

where η ≅ 0.02

T : overvoltage duration TOV : overvoltage value This formula allows the 10 second overvoltage which would cause the same stress on the surge arrester to be calculated for each temporary overvoltage. The duration of the temporary overvoltage must be between several seconds and two to three hours ( U10s = 0.97 × TOV for T = 2 s and U10 s = 114 . × TOV for T = 2 hours ). The rated voltage of the surge arrester will be chosen to be above or equal to the maximum value of the equivalent 10 second voltages: U r ≥ max (U10s ) . We will take

UC ≥

Um 3

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o nominal discharge current I nd

In practice, for the voltage range 1 kV ≤ U m ≤ 52 kV , two values of I nd are available: 5 kA and 10 kA. The value I nd = 10 kA is chosen for areas with a high lightning density.

o discharge class and energy capacity

These are determined by testing or comparison with identical projects.

o mechanical characteristics

The IEC 99-4 and 99-5 standards fix the allowable pressure limit (expressed in "kA") which must be met for the three-phase short circuit at the surge arrester terminals. The surge arrester characteristics will also be checked in relation to: - the ambient temperature - the altitude - the level of pollution - the mechanical resistance to the wind, seismic stress, frost.

o surge arrester protection level

The protection level of the surge arrester at the installation point corresponds to the residual voltage (U rsd ) at its terminals when its nominal discharge current flows through it.

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5.2.3.3.

Installation of HV and MV surge arresters

In HV and MV electrical networks, surge arresters are installed at the entrance to the substation to ensure protection of the substation transformer and equipment. This protection only works if the protection distance and the installation rules are respected.

n protection distance The wave propagation phenomenon studied in § 5.1.5. shows that at the point of reflection (e.g. MV/LV transformer), the overvoltage reaches double the value of the incident wave. The surge arrester peaks at a sparkover voltage U rsd (equal to the residual voltage for ZnO surge arresters). If it is located a considerable distance away, the maximum voltage at the location of the equipment to be protected will thus be 2 U rsd . Now, the equipment impulse withstand is generally lower than 2 U rsd . To overcome this drawback, the surge arrester is installed at a shorter distance away than the "protection" distance D . The surge arrester then undergoes the sum of the incident wave and the reflected wave. It is thus sparked for an incident wave below U rsd . Assuming that at the equipment termination point, the wave is totally reflected, we can show D that the overvoltage in relation to the equipment is limited to U = U rsd + 2 r v r=

dV dt

v

: rise front steepness of the voltage wave, kV/µs : wave propagation speed, m/µs

For a lightning impulse withstand voltage Ul , the surge arrester must therefore be located at a distance D such that: Ursd + 2 r

whence

D ≤ Ul v

U − U rsd D≤ l ⋅v 2r

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Numerical application: Let us consider the example illustrated in figure 5-50: Ul = 125 kV

, case of an MV/LV transformer complying with IEC 76.3

Ursd = 75 kV

, residual voltage of the surge arrester

r = 300 kV / µs

, voltage wave rise front steepness

v = 300 m / µs

, for an overhead line (speed of light)

we then have D ≤

125 − 75 × 300 2 × 300

D ≤ 25 m

The surge arrester must therefore be installed less than 25 m away from the transformer for the overvoltage not to exceed the lightning impulse withstand value.

lightning impulse

A overhead line ZC

D

transformer B

ZC surge arrester

Figure 5-50: protection distance of a surge arrester protecting a transformer fed by an overhead line

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5.2.4.

protection of LV installations

n general LV installations are protected against overvoltages by installing devices in parallel; 3 types of devices are used: - overvoltage limiters located on the secondary of MV/LV transformers (only in an earthing system); they only provide protection against power frequency overvoltages

IT

- low voltage surge arresters installed in LV switchboards or incorporated in loads - surge diverters designed to protect telephone networks, LV terminal boxes and loads.

The main technologies used are: - zener diodes - gas discharge tubes - zinc oxide varistors.

Zener diodes have the drawback of only ensuring the protection of a precise point in the network. The gas discharge tube requires the addition of a varistor to prevent follow current. Variable resistor-type surge arresters are currently the most cost-effective solution owing to their simplicity and reliability

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n LV surge arrester installation rules The equipment is only protected properly if certain installation rules are followed: - rule 1 The length of the connection between the surge arrester and its disconnecting circuit-breaker must be below 0.5 m.

disconnecting circuit-breaker

L < 50 cm

load to be protected

Figure 5-51: diagram of connections

- rule 2 The outgoing feeders of the protected conductors must be connected to the terminals of the surge arrester and its disconnecting circuit-breaker.

- rule 3 The loop surfaces must be reduced by tightly grouping together the incoming, phase, neutral and PE wires.

- rule 4 the incoming wires of the surge arrester (polluted) must be moved away from the protected outgoing wires (healthy) in order to avoid any possible electromagnetic coupling.

- rule 5 The cables must be flattened against the metal structures of the box in order to reduce frame loops and thus benefit from a reduction in disturbances.

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n connection layout according to the earthing system In figures 5-52-a and 5-52-b the connection layouts of the LV surge arrester are shown for different earthing systems.

electrical switchboard

disconnecting circuit-breaker

RCD

equipment to be protected

surge arrester PE

PE

Ph1 Ph2 Ph3 N

LV neutral earth electrode

main earth terminal

load earth electrode

(entrenched loop)

TT earthing system electrical switchboard

disconnecting circuit-breaker equipment to be protected

surge arrester PE PE Ph1 Ph2 Ph3 N PIM

overvoltage limiter

LV neutral earth electrode

main earth terminal

load earth electrode

(entrenched loop)

IT earthing system Figure 5-52-a: connection layout of an LV surge arrester for TT and IT earthing systems

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electrical switchboard

disconnecting circuit-breaker equipment to be protected

surge arrester

PEN Ph1 Ph2 Ph3 PEN

LV neutral earth electrode

main earth terminal

load earth electrode

(entrenched loop)

TNC earthing system

electrical switchboard

disconnecting circuit-breaker equipment to be protected

surge arrester PE PE

Ph1 Ph2 Ph3 N PE LV neutral earth electrode

main earth terminal

(entrenched loop)

load earth electrode

TNS earthing system Figure 5-52-b: connection layout of an LV surge arrester for TNC and TNS earthing systems

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5.3. 5.3.1.

Insulation co-ordination in an industrial electrical network General

Co-ordinating the insulation of an installation consists in determining the insulation characteristics necessary for the various network elements, in view to obtaining a withstand level that matches the normal voltages, as well as the different overvoltages. Its ultimate purpose is to provide dependable and optimised energy distribution. Optimal insulation co-ordination gives the best cost-effective ratio between the different parameters depending on it: - cost of equipment insulation - cost of overvoltage protections - cost of failures (loss of operation and destruction of equipment), taking into account their probability of occurrence.

With the cost of overinsulating equipment being very high, the insulation cannot be rated to withstand the stress of all the overvoltages studied in paragraph 5.1. Overcoming the damaging effects of overvoltages supposes an initial approach which consists in dealing with the phenomena that generate them, which is not always very easy. Indeed, if using the appropriate arc interruption techniques the switchgear switching overvoltages can be limited, it is impossible to prevent lightning strikes.

n clearance (see fig. 5-53) This term covers two notions: - gas clearance (air, SF6, etc.), which is the shortest path between two conductive parts. - creepage distance: this is also the shortest path between two conductors, but following the outer surface of a solid insulating material (e.g. insulator).

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The clearance is directly related to the withstand of the equipment to different overvoltages.

distance in air

creepage distance distance in air

Figure 5-53: air clearance and creepage distance

n overvoltage withstand The overvoltage withstand depends on the type of overvoltage applied (magnitude, wave form, frequency and duration, etc.). It is also influenced by external factors such as: - ageing - environmental conditions (humidity, pollution) - variation in air or insulating gas pressure.

n withstand voltage Electrical equipment is characterised by its withstand voltage to different types of overvoltages. We can thus distinguish: - the power frequency withstand voltage - the switching impulse withstand voltage - the lightning impulse withstand voltage.

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o power frequency withstand voltage

This corresponds to the equipment withstand to power frequency overvoltages likely to occur on the network and the duration of which depends on the network operating and protection mode. The equipment withstand is tested by applying a sinusoidal voltage with a frequency of between 48 Hz and 62 Hz for one minute. The test is valid for nominal network frequencies of 50 Hz and 60 Hz (see IEC 71-1).

o switching impulse withstand voltage

This characterises the equipment withstand to switching impulses (only for equipment with a standard voltage above or equal to 300 kV). The equipment test (see IEC 60-1) is performed by applying a wave with a front time of 250 µs and a time to half-value of 2500 µs.

o lightning impulse withstand voltage

This characterises the equipment withstand to the 1.2 µs / 50 µs lightning voltage wave. This withstand voltage concerns all voltage ranges, including low voltage. o examples of equipment withstand (see table 5-5) Highest voltage for the equipment U m (kV) (1) (r.m.s. value) 3.6

(1)

Standard short-duration power frequency withstand voltage (kV) (r.m.s.)

Standard lightning impulse withstand voltage (kV) (peak value)

10

7.2

20

12

28

17.5

38

24

50

36

70

52 72.5

95 140

20 40 40 60 60 75 95 75 95 95 125 145 145 170 250 325

Um is the highest rms value of the phase-to-phase voltage for which the equipment is specified.

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Table 5-5: standard withstand voltages for 3.6 kV < U m < 72.5 kV

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5.3.2.

Reduction in risks and overvoltage levels

The risks of overvoltages, and consequently the danger they represent for persons and equipment, can be greatly reduced if certain measures of protection are taken: - limiting substation earth electrode resistances in order to reduce power frequency overvoltages - reducing switching overvoltages by choosing suitable switchgear (interruption in SF6) - making lightning impulses flow to earth by a first clipping operation (surge arrester or sparkgap at the entrance to the substation) with limitation of the earth electrode resistances and pylon impedances - limiting the residual voltage from the first clipping operation by HV surge arrester which is transferred to the downstream network by providing a second protection level on the transformer secondary - protection of sensitive equipment in LV (computers, telecommunications, automatic systems, etc.) by connecting series filters and/or overvoltage limiters to it. 5.3.2.1.

Rise in potential of LV exposed conductive parts following an MV fault in the transformer substation

In this paragraph, we propose to study overvoltages in LV caused by an earth fault on the MV side in an MV/LV substation, and the measures to be taken in order to protect equipment and persons, in compliance with IEC 364-4-442. The values of rises in potential of the substation or LV installation exposed conductive parts depend on the values of the earth electrode resistances, the fault current values and the earthing system.

n earthing in transformer substations A single earth electrode must be installed in a transformer substation, to which must be connected: - the transformer tank - the metallic coverings of high voltage cables - the earth conductors of high voltage installations - the exposed conductive parts of high voltage and low voltage equipment - the extraneous conductive parts.

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n symbols In the following paragraphs, the symbols used have the following signification: Im : part of the earth fault current in the high voltage installation which flows through the earth electrode of the transformer substation exposed conductive parts

Re : transformer substation earth electrode resistance V : low voltage installation phase-to-neutral voltage U : low voltage installation phase-to-phase voltage U f : fault voltage in the low voltage installation, between the exposed conductive parts and earth U1 : stress-voltage in the transformer substation low voltage equipment U 2 : stress-voltage in the installation low voltage equipment

n TN − a and IT − a earthing systems (see fig. 5-54) In these two systems, the substation, neutral and installation earth electrodes are the same. Inside the equipotential area, the ground and exposed conductive part potentials increase simultaneously. The touch voltage U f is then zero. On the other hand, outside this area, the ground potential remains equal to that of the remote earth, while the potential of the exposed conductive parts increases to U f = Re I m . Thus, when there are exposed conductive parts outside of the equipotential area and the touch voltage U f = Re I m cannot be cleared in the time defined in tables 2-3-a and 2-3-b, the TN − a and persons.

IT − a

earthing systems are not acceptable in relation to the protection of

To overcome this drawback, the following provisions must be taken: -

TN − a earthing system: the neutral of the LV installation must be connected to a separate earth electrode, which is the case in the TN − b earthing system (see fig. 5-55)

-

IT − a earthing system: the exposed conductive parts of the LV installation must be connected to a separate earth electrode from that of the substation, which is the case in the IT − b earthing system (see fig. 5-56).

TN − b and IT − b earthing systems allow dangerous touch voltages to be cleared but make overvoltages occur: - in the installation LV equipment for the IT − b earthing system - in the substation LV equipment for the TN − b earthing system.

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substation

LV installation

U2

U1

MV

LV

ph 1 ph 2 ph 3

PEN

Uf

0

equipotential zone

Im

U1

Re

Uf

Re I m

Uf

Re I m

outside zone

V

U2

U1 V

TN − a

U2

U1

LV

MV

Z

Uf equipotential zone

Im

Re

outsite zone

U1 V (*) a first LV fault is present

0

U2

3*

U1 V

3*

IT − a

Figure 5-54: rise in potential of TN-a and IT-a earthing systems

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n TN − b , TT − b and IT − c earthing systems (see fig. 5-55) In these three systems, we can see a rise in potential of the exposed conductive parts of the substation U1 such that: U1 = Re I m + V

for TN − b and TT − b earthing systems

U1 = Re Im + V . 3

for IT − c earthing systems with the presence of a first fault on the LV side

Depending on the maximum current value Im , the values of Re must be limited so that U1 remains below the power frequency withstand voltage U tp of the substation equipment. U1 ≤ U tp Table 5-6 gives the maximum values of Re for different values of I m and U tp .

Values at Re not to be exceeded Fault current Im (A)

U tp = 2 000 V

U tp = 4 000 V

U tp = 10 000 V

Class I

Class II

Special class

TN − b ; TT − b ; IT − c

TN − b ; TT − b ; IT − c

TN − b ; TT − b

IT − c

300 A

5.9 Ω

5.3 Ω

12 Ω

30 Ω

1 000 A

1.8 Ω

1.6 Ω

3.6 Ω

10 Ω

5 000 A

0.35 Ω

0.32 Ω

0.72 Ω

2Ω

Table 5-6: maximum values of Re in TN − b , TT − b and IT − c earthing systems

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U1

U2

MV

LV

ph 1 ph 2 ph 3 PEN

Im

RB

Re

U1

Re Im V

U2

V

Uf

0

Uf

TN − b U1

U2

MV

LV

ph 1 ph 2 ph 3 N

Im

Re

RB

U1

Re Im V

U2

V

Uf

0

RA

Uf

TT − b U1

U2

LV

MV

Im

Re

Z

U1

Re I m V 3 *

U2

V 3

Uf

RA If

If

UL

RA U f

(*) a first LV fault is present

IT − c

Figure 5-55: rise in potential in TN − b , TT − b and IT − c earthing systems Publication, traduction et reproduction totales ou partielles de ce document sont rigoureusement interdites sauf autorisation écrite de nos services. The publication, translation and reproduction, either wholly or partly, of this document are not allowed without our written consent.

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n TT − a and IT − b earthing systems In these two cases the substation earth electrode and that of the neutral are common. The LV installation earth electrode is separate. The earth fault current flows through the common earth electrode (neutral/substation). As shown in figure 5-56, we can see that there is a risk of breakdown for the LV equipment whose earth electrode is separate from that of the substation. The following conditions must be met: UtM > Re Im + V

for the TT − a earthing system

and

UtM > Re I m + V 3

for the IT − b earthing system

whence

UtM − V   Re < Im  UtM − V 3  Re <  Im

for the TT − a earthing system for the IT − b earthing system

where: U tM : power frequency withstand voltage of the installation LV equipment equal to 2V + 1000 for V = 220 to 250 V, i.e. 1500 V

Table 5-7 gives the values of Re for different values of Im . TT − a I m = 300 A

IT − b

4Ω

3.5 Ω

I m = 1000 A

1.2 Ω

1Ω

I m = 5000 A

0.24 Ω

0.2 Ω

Table 5-7: maximum values of Re in TT − a and IT − b earthing systems

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substation

LV installation

U1

MV

U2

LV

L1 L2 L3

N

U1 V

Im

U2

R e Im V

Uf

0

Uf

Re

RA

TT − a

U1

U2

LV

MV

Z

U1 V 3

Im

Re

*

U2

Re Im

Uf

RA If

Uf

V 3* If

UL

RA

(*) a first LV fault is present

IT − b

Figure 5-56: Rise in potential in TT − a and IT − b earthing systems

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n recapitulative table of touch voltages and overvoltages which occur for each earthing system

TN − a Touch voltage Overvoltage of LV installation exposed conductive parts Overvoltage of substation exposed conductive parts Y

: yes

N

: no

IT − a

TT − a

IT − b

TN − b

TT − b

IT − c

Y

Y

N

N

N

N

N

N

N

Y

Y

N

N

N

N

N

N

N

Y

Y

Y

Table 5-8: touch voltages and overvoltages which occur for each earthing system

5.3.2.2.

Rise in potential of the LV exposed conductive parts on occurrence of a lightning impulse

When a lightning overvoltage from the distribution network flows to earth in an MV/LV substation through a protection device (surge arrester or MV spark-gap), there follows a rise in potential of the substation LV exposed conductive parts and/or those of the installation which depends on the earthing system. The level of overvoltages transferred in LV depends on the clipped value Ursd and the earth electrode values. To ensure protection of the LV switchgear against these overvoltages, LV surge arresters must be installed and the resistance of the substation earth electrode limited so that the equipment lightning impulse withstand voltage is not exceeded.

n limiting earth electrode impedances As for the case of the MV earth fault, the limit values of the earth electrode impedances are calculated for each earthing system.

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The overvoltage at a point on the network where the impedance changes is given in the relation: v2 =

Z2 2 v1 Z1 + Z2

(see § 5.1.4)

v1 = U rsd : corresponds in this case to the clipped overvoltage

v2 Z1 = Zc Z2 = Ze

: overvoltage of the substation exposed conductive parts : characteristic impedance of the medium voltage line : substation earth electrode impedance

We thus have: v2 =

Ze . 2 Ursd Zc + Ze

The equipment lightning impulse voltage U tc must be above the overvoltage v2 , whence: Ze . 2 Ursd Utc ≥ Zc + Z e Ze ≤

(

Zc 2 Ursd Utc

)

−1

For U rsd = 120 kV

and Z c = 330 Ω , the impulse impedance Z e Z resistance Re measured in low frequency: Re = e . 1.5

is equal to 1.5 times the

The condition on the value of the substation earth electrode impedance is thus: Re ≤

(

Zc

)

U

1.5 × Ursd − 1 tc

The maximum values of Re for the different earthing systems are given in table 5-9. TN − b , TT − b , IT − c

Earthing system

TT − a , IT − b

U tc (kV)

4

8

20

3

Re

3.8

7.7

20.2

2.7

Table 5-9: maximum values of the MV/LV substation earth electrode resistances recommended for limiting MV atmospheric overvoltages transferred in LV

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